9.154. The first design limitation to be considered is related to the heat flux at which a boiling crisis could occur; for a PWR this is the DNB flux, which can be computed by means of a suitable correlation (§9.103). The DNB flux is expected to decrease as the coolant quality increases, since the larger the vapor content of the fluid, the closer it would be to the conditions at which DNB occurs.[15] The quality increases as the enthalpy of the fluid is increased; hence, the enthalpy rise has an important bearing on the value of the DNB flux. In computing this flux for design purposes, the maximum value of the enthalpy, i. e., the value in the hot channel, must therefore be used.

9.155. Since the enthalpy of the coolant increases as it flows upward through the core, the (computed) DNB flux decreases correspondingly, as depicted in Fig. 9.22. Assuming a sinusoidal axial distribution of the vol­umetric heat source, as in §9.146, the heat flux along the hypothetical hot channel will ideally have the same general distribution, as indicated in the figure. The maximum of this curve, representing the highest possible heat flux in the core for normal operating conditions, is equal to the average core heat flux multiplied by the overall heat flux hot-channel factor. In practice, the axial heat flux will not have a sinusoidal distribution, and this must be taken into consideration by the designer. For the present purpose, however, the ideal axial distribution may be assumed.

9.156. Comparison of the computed DNB heat flux with the hot-channel value at any point along the flow channel gives the departure-from-nucleate — boiling ratio (DNBR) or critical heat flux ratio (CHFR) at that point. The results for the whole channel are shown by the uppermost curve in Fig. 9.22. It is seen that the DNBR passes through a minimum. In the interest of reactor safety, an essential requirement of reactor design at present is that the minimum DNBR shall be greater than 1.3 for the hottest channel at the 118 percent overpower specified for a PWR. This is one of the important constraints in reactor core design. It should be noted that the minimum DNBR does not occur at the point of maximum heat flux but more toward the exit where the enthalpy of the coolant is higher and the DNBR flux is lower (§9.103). Factors that change the axial neutron (and heat) flux distribution, e. g., insertion of control rods, will affect the location of the minimum DNBR.

9.157. The design constraint just considered should provide safety from the onset of DNB under normal operating conditions, allowing for a certain amount of overpower that could arise from the instrumental errors or minor reactivity transients. However, no allowance is made for abnormal situa­tions that might result from excessive overpower or inadvertent decrease in the coolant-flow rate. These matters are aspects of reactor safety con­sidered in Chapter 12.

9.158. A somewhat more sophisticated approach preferred by many PWR designers uses the CHF power ratio as a measure of overpower that can be tolerated before the critical heat flux (CHF) condition is reached at the hot channel. In this method, the CHF correlation curve is recalcu­lated by computer, point by point, as the channel power levels under consideration are changed, thus reflecting shifts in mass-flow rate and other parameters. On a thermal flux, coolant enthalpy coordinate system, a family of curves are then plotted showing the proposed channel flux and corresponding variation of CHF, each at the same power level. The CHF power ratio is defined as the power at which the two curves would be tangent to one another, divided by the design or “rated” power. Since relative displacements of the curves with power level are considered, the margin is more physically meaningful than the DNBR [22].

9.159. The central fuel temperature of a cylindrical fuel rod depends on Qa2, where a is the diameter of the fuel pellet, as may be seen from §9.45 et seq. and Example 9.3. By equation (9.20), this is related to the linear heat rate #L, which is a characteristic property of the fuel material for the specified temperature range. Experiments with uranium dioxide, commonly used as fuel in water-cooled (and moderated) reactors, have shown that melting does not occur until the linear heat rate in the fuel exceeds about 70 kW/m. This is, therefore, the maximum permissible linear heat rate anywhere in the reactor core.

9.160. From equation (9.20), qb is equal to Q(ira2) and the heat flux q/A is equal to Qa2/2b, where b is the outer radius of the clad rod. It follows, therefore, that

If the maximum heat flux in a PWR is about 1.5 x 106 W/m2, and the outer diameter of the clad fuel rod is 0.0095 m, i. e., b = 0.00475 m, the corresponding value of qb is found to be about 45 kW/m. It is apparent that the maximum (hot-channel) heat flux in this reactor is moderate. The average linear heat rate, equal to the maximum divided by the heat flux hot-channel factor 2.5, is about 18 kW/m.

9.161. The zirconium alloy (zircaloy) fuel-rod cladding is subject to failure from a variety of causes, including formation of hydrides, stress corrosion, embrittlement, and interaction with the uranium oxide (§7.172). An increase in the cladding temperature appears to result in an enhanced failure rate. There is little danger in normal operation that the cladding will become hot enough to fail, but under accidental conditions, such as a decrease (or cessation) of the coolant flow, the cladding temperature be­comes a critical parameter (Chapter 12).

9.162. Although the temperatures of the fuel and cladding are not likely to represent design constraints, it is nevertheless a common practice to compute the maximum cladding and central fuel temperatures for normal operation and to include them in the list of specifications. The procedure is based on the principles developed in §9.45 et seq. and §9.143 eq seq., with allowance for the variation of the heat conductivity of the fuel material with temperature. The coolant temperature required for the calculations is derived for the hot channel with the minimum flow rate. As seen in §9.149, the maximum cladding surface temperature occurs beyond the point
of maximum heat flux (or input), and so also does the maximum central fuel temperature, although not at the same point as the maximum cladding temperature. Typical maximum values in a PWR are 350°C for the cladding and 1700 to 1900°C for the fuel. The melting point of the uranium dioxide in the fuel pellets is about 2760°C initially, but it drops to some 2650°C during reactor operation as a result of the accumulation of fission products.

Introduction

10.86. The energy derived from the fuel in a nuclear reactor, which is generally sold in the form of electricity, represents a commercial product. The cost of generating this product is made up of several components, one of which is the cost of the fuel. In order to relate fuel costs to other costs, a brief outline will be given of the various contributions to nuclear energy costs. It is customary to apportion the costs of generating electricity, from either nuclear or fossil fuels, to three categories: capital (or construction) costs, expenditures for operation and maintenance, and fuel costs. These will be considered in turn. However, to provide desirable background, we will first review the engineering economics concept of the time value of money.

11.66. The solvent-extraction method for separating the constituents of an aqueous solution can be used when one or more of these constituents are appreciably soluble, whereas the others are much less soluble, in an organic solvent which is essentially immiscible with water. When the or­ganic liquid is brought into intimate contact with the aqueous solution, the substances present will distribute themselves between the organic and aqueous phases. The constituent (or constituents) with the greatest solubility in the organic medium will tend to pass into that phase whereas the others will tend to remain in the aqueous solution. Thus, a partial separation of the constituents of the solution will have been achieved.

11.67. In choosing the organic liquid for a particular solvent-extraction procedure, an important property is its selectivity, that is its ability to extract a particular component (or components) of a solution in preference to all others that are present. The selectivity is expressed by the separation factor, i. e., the ratio of the distribution coefficients of the wanted and unwanted species when equilibrium is attained between the two phases. The distribution coefficient or distribution ratio D is defined as

P _ Cone, of component in organic phase Cone, of component in aqueous phase

at equilibrium, and the separation factor a is given by

_ D (product)

D (impurity) ’

A good solvent for extraction is one for which the distribution coefficient for one component is large and the separation factor is either large or small. In other words, it is desirable that D(product) shall be large, whereas D(impurity) should be small, or vice versa.

11.68. The extraction of an inorganic compound, such as a nitrate, from an aqueous solution by means of an organic solvent is influenced by a number of circumstances. Of particular importance are the presence of (1) salting agents, (2) complex-forming anions, and (3) oxidizing or reducing agents.

11.69. A salting agent is either a salt or an acid, having the same anion as the inorganic compound to be extracted, the presence of which in the aqueous solution increases the distribution ratio. In the extraction of uranyl nitrate, for example, either nitric acid or one if its salts, such as sodium, potassium, calcium, or aluminum nitrate, can serve as a salting agent. These substances are soluble in the aqueous phase but not in the organic solvent.

11.70. The extraction of a specified element from aqueous solution by an organic medium is dependent upon the particular form in which the element is present in the solution. For example, uranyl nitrate hexahydrate can be extracted by certain organic solvents, but the corresponding sulfate is not extractable by these solvents. The addition of a sulfate or other salt of a complex-forming anion to an aqueous solution of uranyl nitrate will thus decrease the extractability of the uranium, since a proportion of the element will be in some form other than the nitrate. The complex-forming anions decrease the distribution coefficient between the organic and aqueous phases, and so their effect is opposite to that of the common-ion salting agents.

11.71. The solvent-extraction processes for separating uranium and plu­tonium from the fission products and then from each other depend on the somewhat unusual chemical behavior of the heavy elements. Starting with actinium (atomic number 89), there is a series of 15 elements, called the actinide series, in the sixth period of the periodic system which resemble the rare-earth (or lanthanide) elements in the fifth period. The lanthanide elements all have similar chemical properties, based on a positive valence of 3, resulting from the presence of three relatively loosely bound outer electrons in each atom. In the analogous actinide series, there are also marked resemblances among the elements, especially in the formation of a tripositive (III) valence state. However, because some of the actinide elements have inner electrons which are not very tightly bound, it is possible to realize tetrapositive (IV), pentapositive (V), and hexapositive (VI) states.

11.72. Although in a given valence state the various actinide elements have similar chemical properties, these properties often are very different in the different oxidation states. For example, the nitrates of the (IV) and (VI) states are appreciably soluble in certain organic liquids, but the nitrates of the (III) states are virtually insoluble in these liquids. The relative stability of the different oxidation states varies with the atomic number of the element. Hence, by the use of appropriate reagents, it is possible to shift the oxidation and reduction states so that two (or more) elements in a given solution will be in different states with differing solubilities in an organic liquid. Separation of the elements can then be accomplished by solvent extraction.

11.73. It is evident from the foregoing discussion that the nitrates of the (IV) and (VI) states of the actinide elements will have large distribution coefficients and hence will be extractable by an organic liquid, but the lower oxidation (III) state will have a smaller distribution coefficient and be less extractable. One consequence of this difference in the distribution coefficients is that after an element has been extracted into an organic medium it can be back-extracted into an aqueous solution if the (IV) or (VI) state is reduced to the (III) state. Suppose two actinide elements have been extracted into an organic solvent in the (IV) or (VI) state. If one of the elements is reduced to the (III) state, it can be separated from the other element by back-extraction into an aqueous solution. The uranium and plutonium in spent reactor fuel are separated from one another in this way.

The purex process

11.74. The “Purex” process, using л-tributyl phosphate (TBP) as the extractant, is typical of solvent-extraction procedures employed in the treat­ment of spent fuel. In the form of nitrates, uranium (VI) and plutonium (IV) can be readily extracted from aqueous solution by TBP, whereas the fission products are taken up to a much smaller extent. TBP is relatively stable in the presence of fairly high concentrations of nitric acid, hence, the latter is used as the salting agent.

11.75. In the first cycle of the Purex process, of which an outline flow sheet is shown in Fig. 11.5, the feed consists of an aqueous solution con­taining uranium (VI), plutonium (IV), and fission product (FP) nitrates plus an excess of nitric acid. Sodium nitrite is added to make sure that the plutonium is entirely in the (IV) state, since this form is best extracted by TBP. The feed solution enters at the middle of the first (extraction) column, while the less dense organic extractant (TBP in a kerosene-type solvent) entering from the bottom flows upward. The uranium (VI) and plutonium (IV) nitrates are thereby extracted from the aqueous solution and pass into the organic medium. In the upper part of the column the organic phase is scrubbed with concentrated nitric acid. Most of the fission products that may have entered the organic solvent are now back-extracted into the aqueous phase, but the nitric acid, which acts as a salting agent, largely prevents back-extraction of the uranium and plutonium. The aqueous ef­fluent (raffinate) from the extraction column contains essentially all the fission products with little or no uranium or plutonium.

11.76. The organic phase containing the uranium and plutonium next passes into the second (partitioning) column where it flows upward and meets the downflowing aqueous strip solution containing a reducing agent to reduce plutonium (IV) to plutonium (III). In a modified Purex process, the reduction is performed electrolytically. The plutonium (III) nitrate is not soluble in the organic medium and so it is back-extracted into the aqueous phase. As this flows downward it is scrubbed with fresh TBP moving upward from the bottom of the column. Any uranium (VI) that has passed into the aqueous solution is thereby returned to the organic phase. The aqueous medium, containing plutonium (III) nitrate, leaves at the bottom of the partition column.

11.77. The organic solution of uranium (VI) nitrate, from which the plutonium and fission products have been almost completely separated, is now transferred to the bottom of the third (stripping) column where it flows upward and is stripped by dilute nitric acid flowing downward. In the absence of a salting agent, the uranium is back-extracted into the aqueous phase and then flows out of the bottom of the column. The spent solvent, leaving at the top, is sent to a recovery plant for purification and subsequent reuse in extraction.

11.78. For further purification, both the aqueous uranium (VI) and plutonium (III) nitrate solutions are submitted to a second and third cycle. The uranium purification in each cycle is essentially identical with the last two stages of the first cycle shown in Fig. 11.5. The aqueous uranium solution is first extracted into the TBP phase and scrubbed with a reducing solution; the organic phase is then stripped by dilute nitric acid in a second column. For each plutonium cycle, the plutonium (III) solution is converted into the (IV) state by means of sodium nitrite and nitric acid, extracted into the TBP medium, and scrubbed with nitric acid in the same column. The organic solution then passes to a stripping column where the plutonium is back-extracted into the aqueous phase by dilute nitric acid. Ion exchange has been employed for the final purification of plutonium following the first cycle of recovery by solvement extraction. The procedure is particu­larly valuable for the separation of zirconium and ruthenium, and their daughter products niobium and rhodium, respectively. The common prac­tice, however, is to use two additional stages of solvent extraction to remove residual fission products, as stated above.

12.91. The temperature of the fuel rod cladding would rise to a maxi­mum in the reflood stage after an LOCA. The integrity of the cladding must be maintained so that it would not fragment under the thermal stress subsequently imposed when the very hot fuel rods are quenched. Embrit­tlement of zircaloy, which could lead to fragmentation, is a function of the temperature and degree of oxidation. Criteria (1) and (2) are intended to preclude the embrittlement (and melting) of the cladding.

12.92. The objective of criterion (3), which limits the amount of hy­drogen gas produced by the zirconium-steam (or — water) reaction, is to keep the concentration of the gas in the containment vessel well below that at which a hydrogen-air mixture would ignite.

12.93. Local swelling (or ballooning) of the cladding, which might result from the expansion of contained fission-product gases, could affect the flow of coolant water through the core. According to criterion (4), cal­culations must show that despite changes in its internal geometry, the core should remain coolable during the reflood stage. The purpose of criterion (5) is obvious and does not require discussion.

12.94. The conservative nature of the foregoing criteria assures an ad­equate margin of performance of the ECCS should a design basis LOCA

occur. This margin is provided by the criteria themselves as well as by conservative features of the evaluation models (§12.132 et seq.). The con­servative features include the following:

1. The calculation of stored heat in the fuel is based on a preaccident reactor power level of 102 percent of the maximum operating power, with the highest allowed peaking factor (§9.167) and the lowest estimated thermal conductance between the fuel pellets and the cladding (§9.50).

2. Heat transfer during blowdown must be calculated using NRC-approved realistic models having an extensive experimental basis.

3. The fission-product decay heat is taken to be 20 percent greater than in the ANS standard (§2.217). The use of the Baker-Just relationship (§7.130) for calculating the heat generation rate from the zirconium-water reaction is also conservative.

4. The peak cladding temperature of 2200°F (1204°C) refers to the hottest region of the hottest fuel rod. This criterion ensures that there would be very little damage to the reactore core.

12.187. The destruction of Unit 4 of the Chernobyl Atomic Power Sta­tion in Ukraine on April 26, 1986 was the worst nightmare of nuclear engineers throughout the world. Attempts to convince a skeptical public that nuclear power is indeed a safe energy option would now be more difficult than ever, even though the reactor design is unique to the former Soviet Union and inherently unstable. A first step in discussing the accident and its impact is to become familiar with the system.

12.188. The reactor, one of the RBMK-1000 type, was boiling light — water-cooled and graphite-moderated. Fuel element subassemblies, within pressure tubes through which the coolant flowed, contained 18 rods ar­ranged in concentric rings. At full power, the coolant channel exit steam quality was 14 percent. The core, 7 m high and 11.8 m in diameter, con­tained 1661 such channels. Refueling was accomplished at full power by a machine in a bay above the core, as shown in the Fig. 12.14 plant arrange­ment. A 2 percent enrichment was used, with 211 control rods of various types needed for reactivity compensation. The rated thermal power was 3200 MW.

12.189. As a result of the massive size of the reactor and fuel handling equipment, it was considered impractical to have a complete containment such as used for all U. S. reactors. Instead, an accident localization system, as indicated in Fig. 12.14, was provided. This consisted of various sealed compartments that enclose the circulating pumps, large piping, and other components, the failure of which could lead to a LOCA. This compartment system vents to a suppression pool. The cooling channel refueling con­nections were unprotected. Hence, the accident localization system was designed for a different type of accident, i. e., a large pipe LOCA, rather than the massive fuel channel failure that occurred.

12.190. Since graphite is the moderator in the RBMK-1000 reactor, voiding of the coolant, which is a neutron absorber, increases the reactivity. This positive reactivity coefficient was significantly greater at low power. Therefore, 700 MW(t) was specified as the minimum permissible contin­uous operating power level. An interesting positive reactivity effect is in­troduced when control rods are inserted as a result of the displacement of water at the bottom of the control rod channels by graphite control rod

Primary coolant boundary areas not protected by AtS Semi protected mm

Protected areas

followers. During post-accident analysis, it was also concluded that the emergency rod insertion rate was slow by western standards.

12.191. The accident occurred during a test of turbine-generator coast­ing-down power, which was used to drive an emergency feedwater pump for about 1 minute in the event of the loss of off-site power, with some of the desired electrical conditions simulated. Although the intended reactor power for the test was 700 MW(th), errors by the operators resulted in a power loss to 30 MW(th), where xenon growth, particularly at the bottom of the core, from the previous higher power operation made it difficult to increase the power without withdrawing almost all of the control rods. Even with such action, only a power level of 200 MW(th) could be achieved, a level in violation of operating procedures because of the inherent insta­bility of the reactor.

12.192. The test plan was initiated at this low power level by starting additional recirculating pumps as called for by an electrical simulation, which had the effect of reducing core voids and causing additional control rod withdrawal. As a result of operating difficulties, various protection devices were blocked out by the operators. When the recirculating pumps were allowed to coast down, as planned, coolant flow decreased and voids re-formed very rapidly in the pressure tubes, which increased reactivity because of the positive void coefficient, particularly at the bottom of the core. An emergency scram (trip) was initiated manually almost immedi­ately, but the almost fully withdrawn rods could not be inserted fast enough to prevent a prompt critical power excursion. In fact, the rod insertion introduced some additional reactivity at the core bottom as a result of water displacement by the graphite followers.

12.193. The rapid vaporization of the coolant in the pressure tubes generated a shock wave that ruptured most of the tubes. Apparently, there were two excursions, seconds apart, the second a result of almost complete coolant voiding. The fuel became molten and generated an immense quan­tity of steam, which blew the 1000-ton steel and concrete biological shield off the top of the reactor. Hydrogen, formed by the reaction of fragmented cladding and water, exploded, severely damaging the building. Pressure tube ruptures provided an inlet and an outlet for air to feed combustion of the graphite, which was probably ignited by the exothermic zirconium — niobium oxidation reaction. The fire continued for several days and cer­tainly complicated the management of the accident. The accident was cat­egorized as level 7 on the International Nuclear Event Scale (INES).[25]

12.194. In late 1986, after radiation levels had decreased somewhat, the damaged reactor building was enclosed by a concrete and steel shell, or “sarcophagus.” Several years later, explorations inside the sarcophagus revealed that about 96 percent of the original fuel is contained in solidified “lava flows” in chambers below the reactor vault and in the form of dust and particles distributed inside the building. Since it was necessary to construct the sarcophagus as an emergency measure upon damaged foun­dations, it is expected that replacement will eventually be necessary.

12.195. The total release of particulate fuel from the core is estimated at 3.5 percent of the original inventory. This corresponds to a radioactivity release into the environment on the order of 50 million curies. Fallout over parts of the former Soviet Union and other countries was widespread. Considering the most biologically sensitive fission products, 100 percent of the rare gases was released, about 20 percent of the iodine, and roughly 13 percent of the cesium. The estimated 2 million curie release of the 30- year half-life cesium-137 is the most significant long-term contamination contribution. The initial 32 fatalities all occurred on-site. The long-term consequences of the exposure received by about 200 plant workers treated for radiation sickness at the time of the accident and off-site exposure to fallout by some of the nearby general population remains uncertain.

8.21. As part of our study of reactor control, we learned how feedback could affect the behavior of a system, particularly its stability (§5.132).

L

Should the output of our system serve as the input to a second system or subsystem and then have one of the second system outputs return as input to the original system, we have a feedback effect. In the case of reactor control, the second system could either reinforce or inhibit an original disturbance signal, leading to an unstable or a stable condition, respectively.

8.22. In the reactor design logical flow diagram shown in Fig. 8.1, the required fuel loading, flux pattern, and power distribution evolve from the nuclear design, and some feedback is provided to both the nuclear data and thermal-hydraulic design activities. Thus, we have a typical procedure initiated by a trial design which is then refined by feedback. Although not shown in the figure, the process involves an evaluative stage in which results are compared with goals. This procedure lends itself to automation using standard computer programming practice.

9.87. The contribution of high molecular conductivity results in heat transfer coefficients for a sodium-cooled core rod bundle that may be larger than those for a pressurized water reactor by a factor of 2 or 3. Uncertainty in prediction methods can be tolerated by the designer since such high values result in low temperature differences between the heat transfer surface and the bulk of the flowing fluid. A number of semiempirical correlations of experimental data are described in the literature [9].

9.88. Sodium-cooled fast reactors have a power density several times larger than that for a PWR. Also, sodium boiling in the core increases reactivity and must be avoided (§5.129). Hence, the designer needs to consider a number of heat-transfer problems that are only indirectly related to the heat-transfer coefficient based on average conditions. Attention must be given to local effects around a rod’s periphery, local flow restrictions, and a variety of safety-related questions arising from the possibility of accidental loss of coolant flow.

10.15. The useful lifetime of fuel in a reactor is measured by the burnup, expressed as the total amount of thermal energy generated per unit quantity of heavy element charged to the core. In LWRs, the freshly loaded fuel consists of plutonium-free uranium oxide. Hence, the burnup is expressed as the thermal energy produced during the lifetime of the fuel per unit mass of uranium initially charged to the reactor even though significant plutonium is produced during this time. Alternative methods for expressing burnup, especially in irradiation effects studies, are (1) number of fissions per m3 of fuel and (2) percentage of heavy-element atoms that have under­gone fission. As seen in Example 1.2, for a reactor with uranium dioxide fuel, e. g., a LWR, a burnup of 2.6 TJ/kg U (or 30,000 MW • d/t) is equivalent to about 7.4 x 1026 fissions/m3 of fuel and to about 3 percent of atoms fissioned as uranium-235 plus plutonium-239 and plutonium-241.

10.16. There are several reasons why reactor fuel must be discharged long before the fissile and fertile materials are consumed. In the first place, accumulation of fission products and of the isotopes of heavy elements, which act as neutron poisons, and depletion of the fissile species, e. g., uranium-235 and plutonium-239, can decrease the reactivity to such an extent that the operational requirements of the reactor will no longer be satisfied. Furthermore, materials performance levels may deteriorate with continued exposure so that they become unacceptable. For example, leak­age of fission products into the coolant is not tolerated. The relevant fuel performance considerations are described in §7.170 et seq.

10.17. Generally, a high burnup is desirable so that unit costs associated with the preparation of the fuel could be spread over more units of energy generated. When fuel reprocessing was considered, cost studies showed that the increased burnup benefits leveled at a burnup in the vicinity of 2.6 TJ/kg as a result of the influence of enrichment charges [2]. However, with present “once-through” (no reprocessing) practice in the United States there is an incentive to extend the burnup in LWRs until the fuel materials performance levels are no longer acceptable. On the other hand, as burnup is extended, the higher enrichment needed makes the in-core fuel man­agement design more difficult (§10.26 et seq.). In addition to more efficient resource utilization, extended burnup reduces the number of discharged fuel assemblies that must be stored per amount of generated energy. Ex­perimental burnups on the order of 4.3 TJ/kg have been obtained in some LWRs [3].

10.116. In designing a specific operation, attention is given to the above considerations. For example, a high-neutron-leakage geometry can be pro­vided. Neutron absorbers, such as cadmium or boron, either in solution or in a geometric array, are useful. Accumulations of fissile material that might cause problems can be avoided by careful administrative control.

10.117. Design and analysis go hand-in-hand. A prediction of criticality is necessary to confirm the adequacy of design options. This is normally done by machine calculations using suitable codes supplemented by data from critical experiments. Since a nuclear criticality safety calculation should predict a noncritical condition with high reliability, conservative values of constants and inputs should be selected when uncertainty exists. However, since safety margins are provided in the design, the calculation methods can be relatively unsophisticated, provided that the confidence level of the results is known.

1.

INTRODUCTION

12.9. The first goal in reactor safety is to prevent accidents from oc­curring. This goal has two aspects. First, the reactor system needs to be designed, constructed, and operated so that the chances of a malfunction or operational error are very small. Since some equipment failures and operational mistakes are inevitable during the service lifetime of such a complex system as a nuclear power plant, the second aspect of the pre­vention goal is to provide “self-healing” features that will cope with such incidents.

12.10. The attainment of a reliable system by conservative design re­quires anticipating possible modes of failure and making provision for them, meeting demanding quality standards, and adhering to applicable regulatory requirements. Details of system design are beyond the scope of this treatment. However, we will examine some accident scenarios, the prevention of which must be demonstrated as part of the licensing pro­cedure. The need for quality assurance is also enforced as a regulatory requirement.